Nickel-based alloy

Abstract

4.0 to 5.6 mass% aluminum, 0.0 to 1.0 mass% titanium, 0.0 to 4.0 mass% niobium, 0.0 to 11.9 mass% tantalum, 2.0 to 12.7 mass% tungsten, 0.0 to 3.0 mass% molybdenum, 0.0 to 22.0 mass% cobalt, 6.0 to 16.7 mass% chromium, 0.02 to 0.35 mass% carbon, 0.001 to 0.2 mass% boron, 0.00 to 0.5 mass% zirconium, 0.0 to 3.0 mass% rhenium, 0.0 to 2.0 mass% ruthenium, 0.0 to 3.0 mass% iridium, 0.0 to 0.5 mass% vanadium, 0.0 to 1.0 mass% palladium, 0.0 to 1.0 mass% platinum, 0.0 to 0.5 mass% silicon, 0.0 to 0.1 mass% yttrium, 0.0 to 0.1 mass% lanthanum, 0.0 to 0.1 mass% cerium, 0.0 to 0.003 mass% sulfur, 0.0 to 0.25 mass% manganese, 0.0 to 0.1 mass% magnesium, 0.0 to 4.0 mass% iron, 0.0 to 0.5 mass% copper, 0.0 to 1.0 mass% hafnium, and the balance being nickel and unavoidable impurities, and the mass% of niobium and tantalum contained in the alloy are W, respectively. Nb , W Ta A nickel-based alloy composition that satisfies the following formula: 1.1≦0.3W Nb +0.15W Ta

Description

[Technical Field]

[0001] The present invention relates to nickel-based superalloy compositions designed for application in additive manufacturing (AM) processes. Examples of such processes include, but are not limited to, powder bed-based AM methods (e.g., selective laser melting, electron beam melting) and directed metal deposition methods (e.g., powder deposition and wire-based methods). [Background technology]

[0002] Currently, there is a trend to transfer nickel-based superalloys, successfully manufactured by casting or forging, to additive manufacturing (AM) processes. However, this has proven to be largely inappropriate. This is because many of the material properties necessary to facilitate machining in AM processes are not met by the aforementioned alloys, making machining significantly more difficult and resulting in materials that lack the expected structural integrity.

[0003] In particular, developing alloys with a high γ' volume fraction for additive manufacturing processes has been a major challenge, as these alloys are often classified as "non-weldable." Typically, these alloys are processed using investment casting. Table 1 lists examples of common alloys used in investment casting processes. Table 1 shows the nominal composition in mass percent of conventional high γ' volume fraction alloys.

[0004] Extensive research has been conducted on applying the alloys listed in Table 1 to additive manufacturing. Although these alloys may be "non-weldable" and difficult to process, it has been shown that certain defect mechanisms can be limited. For example, while these alloys are susceptible to strain aging cracking, careful control of AM conditions (scan plan, heat input, etc.) and post-processing conditions (heat treatment window and heating rate) during processing can minimize the risk of strain aging cracking to some extent. Similarly, resistance to strain aging cracking can be improved by designing the part geometry to reduce the effects of stress concentration features such as notches.

[0005] Strain aging cracking is caused by two main factors: residual strain and γ' precipitation. The magnitude of residual strain can be easily reduced by preheating the powder bed, which minimizes the temperature range of thermal shrinkage. Illston (Patent No. US9352421B2) has shown that process control (particularly the use of a thin powder bed and intentionally overlapping laser scans) can minimize the accumulation of residual strain, thus improving the printability of high γ' superalloys. Etter et al (Patent No. 9670572B2) have also shown that the risk of strain aging cracking can be reduced by rapidly increasing the temperature to the stress relaxation heat treatment temperature after AM.

[0006] Another mechanism that cannot be easily reduced by process optimization is hot cracking. Hot cracking occurs in the final stage of the solidification process and is highly dependent on the chemical properties of the alloy. [Overview of the project] [Problems that the invention aims to solve]

[0007] The object of the present invention is to develop a high γ' volume fraction alloy that is particularly resistant to the mechanism of hot cracking by tuning the chemical properties of the alloy to overcome the undesirable damage mechanism of hot cracking.

[0008] Combined with this improved workability, the alloy of the present invention preferably has a very high level of oxidation resistance (achieved by having a sufficient level of aluminum to form a protective aluminum-based oxide scale) and a high level of creep resistance.

[0009] [Table 1] [Means for solving the problem]

[0010] According to the present invention, aluminum of 4.0 to 5.6% by mass, titanium of 0.0 to 1.0% by mass, niobium of 0.0 to 4.0% by mass, tantalum of 0.0 to 11.9% by mass, tungsten of 2.0 to 12.7% by mass, molybdenum of 0.0 to 3.0% by mass, cobalt of 0.0 to 22.0% by mass, chromium of 6.0 to 16.7% by mass, carbon of 0.02 to 0.35% by mass, boron of 0.001 to 0.2% by mass, zirconium of 0.00 to 0.5% by mass, rhenium of 0.0 to 3.0% by mass, ruthenium of 0.0 to 2.0% by mass, iridium of 0.0 to 3.0% by mass, vanadium of 0.0 to 0.5% by mass, palladium of 0.0 to 1.0% by mass, platinum of 0.0 to 1.0% by mass, silicon of 0.0 to 0.5% by mass, yttrium of 0.0 to 0.1% by mass, lanthanum of 0.0 to 0.1% by mass, cerium of 0.0 to 0.1% by mass, sulfur of 0.0 to 0.003% by mass, manganese of 0.0 to 0.25% by mass, magnesium of 0.0 to 0.1% by mass, iron of 0.0 to 4.0% by mass, copper of 0.0 to 0.5% by mass, and hafnium of 0.0 to 1.0% by mass are included, with the balance being nickel and unavoidable impurities, and the mass percentages of niobium and tantalum contained in the alloy being W Nb 、W Ta respectively. Then, a nickel-based alloy composition satisfying the following formula is provided. 1.1 ≦ 0.3W Nb + 0.15W Ta Such a nickel-based alloy has excellent hot cracking resistance, along with appropriate creep resistance, appropriate strength, appropriate strain age cracking resistance, appropriate microstructure stability, an acceptable solidification temperature range, an acceptable density, and an acceptable cost.

[0011] In one embodiment, when the mass percentages of niobium and tantalum contained in the alloy are W Nb 、W Ta respectively, the following formula is satisfied. 1.15 ≦ 0.3W Nb + 0.15W Ta Preferably, the following formula is satisfied. 1.2 ≦ 0.3W Nb + 0.15W Ta More preferably, the following equation is satisfied. 1.4 ≤ 0.3W Nb +0.15W Ta More preferably, the following equation is satisfied. 1.8 ≤ 0.3W Nb +0.15W Ta Such alloys exhibit even greater resistance to thermal cracking.

[0012] In one embodiment, the mass percentages of aluminum, titanium, niobium, and tantalum contained in the alloy are each set to W Al , W Ti , W Nb and W Ta Therefore, the following equation is satisfied. 5.6≦W Al +0.5W Ti +0.3W Nb +0.15W Ta ≤7.0 Preferably, the following equation is satisfied. 5.6 ≤ W Al +0.5W Ti +0.3W Nb +0.15W Ta ≤6.5 Such alloys exhibit improved resistance to strain-induced cracking.

[0013] In one embodiment, the mass percentages of tungsten and molybdenum contained in the alloy are W, respectively. W , W Mo Therefore, the following equation is satisfied. W W +0.65W Mo ≥4.0 Preferably, the following equation is satisfied. W W +0.65W Mo ≥6.0 More preferably, the following equation is satisfied. W W +0.65W Mo ≥8.0 Such alloys exhibit improved creep resistance.

[0014] In one embodiment, the chromium content of the nickel-based alloy composition is 8.0% or more by mass, preferably 10.0% or more. Such alloys exhibit improved oxidation resistance and corrosion resistance.

[0015] In one embodiment, the chromium content of the nickel-based alloy composition is 14.7% or less by mass, preferably 13.8% or less. Such alloys exhibit improved microstructural stability.

[0016] In one embodiment, the tantalum content of the nickel-based alloy composition is 9.8% or less by mass, preferably 9.2% or less, and more preferably 7.1% or less. Such alloys have low density, and the tungsten content can be optionally increased. Increasing the tungsten content helps to improve creep resistance.

[0017] In one embodiment, the molybdenum content of the nickel-based alloy composition is 2.0% or less by mass, preferably 1.8% or less. Such alloys exhibit improved heat cracking resistance.

[0018] In one embodiment, the titanium content of the nickel-based alloy composition is 0.5% or less by mass, preferably 0.1% or less. Such an alloy has better oxidation resistance.

[0019] In one embodiment, the tungsten content of the nickel-based alloy composition is 10.7% or less by mass, preferably 8.7% or less. Such alloys exhibit improved microstructural stability.

[0020] In one embodiment, the niobium content of the nickel-based alloy composition is 3.0% or less by mass. Such an alloy exhibits improved oxidation resistance.

[0021] In one embodiment, the nickel-based alloy composition contains one or both platinum and palladium in a mass percentage of 0.5% or less. Such alloys result in even lower costs.

[0022] In one embodiment, the nickel-based alloy composition contains 4.5% or more aluminum by mass. Such an alloy exhibits improved corrosion resistance.

[0023] In one embodiment, the aluminum content of the nickel-based alloy composition is 5.3% or less by mass, preferably 5.1% or less, and more preferably 4.6% or less. Such alloys exhibit improved resistance to strain aging cracking, in addition to improved thermal cracking resistance.

[0024] In one embodiment, the cobalt content of the nickel-based alloy composition is 15.0% or less by mass. In such an alloy, the solidification temperature range decreases.

[0025] In one embodiment, the tantalum content of the nickel-based alloy composition is 1.1% or more by mass, preferably 3.7% or more, and more preferably 5.8% or more. In such alloys, the amount of niobium can be reduced without decreasing the thermal cracking resistance, or thermal cracking resistance can be further improved by combining it with a higher level of niobium.

[0026] In one embodiment, the tungsten content of the nickel-based alloy composition is 2.7% or more by mass, preferably 2.8% or more, more preferably 4.7% or more, even more preferably 5.1% or more, and most preferably 5.2% or more. Such an alloy exhibits improved creep resistance.

[0027] In one embodiment, the mass percentages of tantalum and tungsten contained in the alloy are W, respectively. Ta and W W Then, the following equation is satisfied W W +W Ta ≤13.9 Preferably, the following equation is satisfied. W W +W Ta ≤11.8 These alloys have a lower density.

[0028] In one embodiment, the mass percentages of chromium, molybdenum, and tungsten contained in the alloy are W, respectively. Cr , W Mo and W W Therefore, the following equation is satisfied. W W +W Cr +1.7W Mo ≤18.7 Preferably, the following equation is satisfied. W W +W Cr +1.7W Mo ≤17.8 In such alloys, microstructural stability is improved. [Brief explanation of the drawing]

[0029] [Figure 1] Figure 1 shows the calculated strain-aging merit index and hot cracking index for several commercially used superalloys (including the alloys listed in Table 1). Figure 1 also shows the limits of creep resistance, strain-aging cracking, and hot cracking. The target region in this invention is indicated by hatching. [Figure 2] Figure 2 is an contour plot showing the influence of aluminum, niobium (following the relationship 0.3WNb + 0.15WTa), and tantalum, which are γ'-forming elements, on the γ' volume fraction. This contour plot was obtained by phase equilibrium calculations performed at 900°C, with the titanium content fixed at 0.0 mass% in the alloy within the alloy design domain. The graph shows different limitations on the strain aging index and hot cracking index. [Figure 3] Figure 3 is an contour plot showing the influence of γ'-forming elements, namely aluminum, niobium (following the relationship 0.3WNb + 0.15WTa), and tantalum, on the hot cracking index. This contour plot was obtained by phase equilibrium calculations performed at 900°C, with the titanium content fixed at 0.0 mass% in alloys within the alloy design domain. [Figure 4]Figure 4 is an contour plot showing the effects of γ' volume fraction and creep merit index on creep temperature capability. These contours are normalized to IN713C and show the predicted positions for the alloys listed in Table 1. [Figure 5] Figure 5 is an contour plot showing the effects of molybdenum and tungsten on the creep merit index. This graph illustrates the stability limits at various concentrations of chromium. [Figure 6] Figure 6 is an contour plot showing the effects of molybdenum and tungsten on alloy stability (in terms of Md number) when the chromium content is fixed at 6.0 mass%. This graph shows various levels of creep merit index. [Figure 7] Figure 7 is an contour plot showing the effects of molybdenum and tungsten on alloy stability (in terms of Md number) when the chromium content is fixed at 8.0 mass%. This graph shows various levels of creep merit index. [Figure 8] Figure 8 is an contour plot showing the effects of molybdenum and tungsten on alloy stability (in terms of Md number) when the chromium content is fixed at 10.0 mass%. This graph shows various levels of creep merit index. [Figure 9] Figure 9 is an contour plot showing the effects of molybdenum and tungsten on alloy stability (in terms of Md number) when the chromium content is fixed at 12.0 mass%. This graph shows various levels of creep merit index. [Figure 10] Figure 10 is an contour plot showing the effects of molybdenum and tungsten on alloy stability (in terms of Md number) when the chromium content is fixed at 14.0 mass%. This graph shows various levels of creep merit index. [Figure 11]Figure 11 is an contour plot showing the effects of molybdenum and tungsten on alloy stability (in terms of Md number) when the chromium content is fixed at 16.0 mass%. This graph shows various levels of creep merit index. [Figure 12] Figure 12 is an contour plot showing the effects of tantalum and tungsten on alloy density. [Figure 13] Figure 13 is an contour plot showing the effects of the strain aging cracking index and cobalt on the solidification range of the alloy. [Figure 14] Figure 14 shows a micrograph of a nickel-based superalloy manufactured using a powder bed-based AM method. The influence of the hot cracking index (HCI) on susceptibility to cracking during AM treatment was clarified, which rationalizes the preferred limitation of the hot cracking index (1.5 or less) in the present invention. [Modes for carrying out the invention]

[0030] Traditionally, nickel-based superalloys have been designed based on empiricism. Therefore, the chemical compositions of nickel-based superalloys have been determined through time-consuming and expensive experimental development, involving small-scale processing of limited amounts of material and subsequent characterization of its behavior. Subsequently, the alloy composition found to exhibit the best, i.e., most desirable combination of properties, is adopted. The existence of numerous alloying elements capable of achieving this combination suggests that these alloys are not fully optimized and that more improved alloys likely exist.

[0031] In superalloys, chromium (Cr) and aluminum (Al) are generally added to impart oxidation / corrosion resistance, and cobalt (Co) is added to improve resistance to sulfidation. Molybdenum (Mo), tungsten (W), and cobalt are introduced for creep resistance because these elements inhibit the thermally activated processes (e.g., dislocation elevation) that determine the rate of creep deformation. Aluminum (Al), tantalum (Ta), niobium (Nb), and titanium (Ti) are introduced to increase static and cyclic strength because these elements promote the formation of the precipitation-hardened gamma prime (γ'). This precipitation phase is coherent with the face-centered cubic (FCC) matrix phase called gamma (γ).

[0032] In this specification, a model-based method used to identify new grades of nickel-based superalloys is referred to as the “Alloy Design” (ABD) method. This method utilizes a computational materials model framework for estimating design-related properties across a very broad range of compositions. In principle, this alloy design tool enables the solution of so-called inverse problems; that is, it allows for the identification of the optimal alloy composition that best satisfies specified design constraints.

[0033] The first step in the design process is to define the elemental table and the upper and lower limits of the compositional constraints associated with that table. In this invention, the elemental composition constraints for each element when adding each element, called the "alloy design region," are considered. These compositional constraints are described in detail in Table 2. Table 2 shows the alloy design region in mass percent, as investigated using the "alloy design" method.

[0034] [Table 2]

[0035] The remainder is nickel. The levels of carbon, boron, and zirconium were fixed at 0.06%, 0.015%, and 0.06%, respectively.

[0036] The second step involves thermodynamic calculations to determine the phase diagram and thermodynamic properties of a specific alloy composition. This is often referred to as the CALPHAD method (Calculation of PHAse Diagrams). Performing these calculations at the typical operating temperature of the new alloy (900°C) provides information about the phase equilibrium (microstructure).

[0037] The third step involves identifying alloy compositions with the desired microstructure. For nickel-based superalloys requiring excellent resistance to creep deformation, the creep rupture lifetime gradually improves as the volume fraction of the precipitation-hardening phase γ' increases. The range of γ' volume fraction that provides the most beneficial creep rupture lifetime is 60–70%. When the γ' volume fraction exceeds 70%, a decrease in creep resistance is observed.

[0038] Furthermore, the γ / γ' lattice inconsistency must follow the smaller of either positive or negative values, as it would result in a loss of coherence. Therefore, the restriction depends on the absolute value of that value. The lattice inconsistency δ is defined as the mismatch between the γ phase and the γ' phase and is calculated by the following formula.

[0039]

number

[0040] Here, α γ and α γ´ These are the lattice constants of the γ phase and the γ' phase.

[0041] Alloys based on unsuitable microstructures are also rejected by estimated susceptibility to morphologically close-packed (TCP) phases. Using CALPHAD modeling in this calculation, the formation of harmful TCP phases sigma (σ), Ρ, and mu (μ) is predicted.

[0042] Therefore, this model identifies all compositions within the design domain for which the calculated volume fraction of γ' is a desired value. In these compositions, the lattice imperfection of γ' is less than a predetermined absolute value, and the total volume fraction of the TCP phase is less than a predetermined magnitude.

[0043] In the fourth stage, merit indices are estimated for the identified alloy compositions remaining in the dataset. Examples of merit indices include the creep merit index (indicating the creep resistance of the alloy based on average composition only), the strength merit index (indicating the precipitation yield strength of the alloy based on average composition only), the solid solution merit index (indicating the solid solution yield strength of the alloy based on average composition only), density, and cost.

[0044] In the fifth stage, the calculated merit index is compared to the constraints on the desired behavior, and these design constraints are considered boundary conditions for the problem. All compositions that do not satisfy the boundary conditions are rejected. At this stage, the size of the test dataset becomes very small.

[0045] The final, sixth stage involves analyzing the dataset of remaining compositions. This analysis can be carried out in various ways. One method is to classify the alloys via a database based on which alloys exhibit the highest merit index. Alloys exhibiting the highest merit index are, for example, the lightest alloys, the alloys with the highest creep resistance, the alloys with the highest oxidation resistance, and the cheapest alloys. Alternatively, the database may be used to determine the relative trade-offs in performance resulting from different combinations of properties.

[0046] I will explain seven examples of the merit index.

[0047] The first merit index is the creep merit index. The most important observation is that the time-dependent deformation (i.e., creep) of nickel-based superalloys is caused by dislocation creep associated with initial activity limited to the γ phase. Therefore, as the proportion of the γ' phase increases, dislocation segments are rapidly fixed at the γ / γ' interface. The rate-determining step is the detachment of the trapped dislocation configuration from the γ / γ' interface. This depends on the local chemistry (in this case, the composition of the γ phase) which has a significant influence on the creep properties of the alloy composition.

[0048] A physics-based microstructure model is one in which the load is uniaxial. <001> When aligned with the crystallographic direction, creep strain ε · This is used to calculate the accumulation rate. The set equation is as follows:

[0049]

number

[0050] Here ρ m φ is the mobile dislocation density. p σ is the volume fraction of the γ' phase, and ω is the width of the matrix channel. Terms σ and Τ are the applied stress and temperature, respectively. Terms b and k are the Burgers vector and Boltzmann constant, respectively. Term K CF This is the constraint coefficient.

[0051]

number

[0052] term K CF This indicates the proximity of cubic particles within these alloys. Equation 3 shows the dislocation multiplication process, which requires estimation of the multiplication parameter C and the initial dislocation density. Term D eff This is the effective diffusion rate that controls the upward movement process at the particle / matrix interface.

[0053] Furthermore, in the above description, the composition dependence is expressed by two terms φ p and D effIt arises from this. Therefore, assuming that the microstructure is constant (most of the microstructure is controlled by heat treatment), φ p Since it is fixed, the dependence on the chemical composition is D eff This is caused by... It turns out that for the purposes of the alloy design modeling described here, it is not necessary to perform the full integration of Equations 2 and 3 for each prototype alloy composition. Instead, the first merit index M, which needs to be maximized, is... creep M is used. creep It can be calculated using the following formula.

[0054]

number

[0055] Here, x i This is the atomic fraction of solute i in the γ phase. i ~ This is the appropriate interdiffusion coefficient.

[0056] The second merit index is the strength merit index. In the case of high-nickel-based superalloys, the majority of the strength comes from the precipitated phase. Therefore, optimizing the alloy composition to maximize the precipitation strength is an important design consideration. Based on hardening theory, the strength merit index M strength The following is proposed. This index takes into account the maximum possible precipitation strength (determined as the point at which dislocation shear transitions from weak bonds to strong bonds occur) and is approximated by the following formula.

[0057]

number

[0058] Here, M- is the Taylor coefficient, γ APB φ is the anti-phase boundary (APB) energy. p is the volume fraction of the γ' phase, and b is the Burgers vector.

[0059] From equation (5), it is clear that the defect energy in the γ' phase (e.g., the anti-phase boundary APB energy) has a significant impact on the deformation behavior of nickel-based superalloys. Increasing the APB energy has been found to improve mechanical properties, including tensile strength and resistance to creep deformation. Studies of APB energy have been conducted for many Ni-Al-X systems using density functional theory. This study calculated the effect of ternary elements on the APB energy of the γ' phase, and assumed a linear superposition of the effects of each ternary element addition when considering complex multi-component systems. As a result, the following equation was derived.

[0060]

number

[0061] Here, x Cr , x Mo , x W , x Ta , x Nb and x Ti These represent the atomic percentage concentrations of chromium, molybdenum, tungsten, tantalum, niobium, and titanium in the γ' phase, respectively. The composition in the γ' phase is determined by phase equilibrium calculations.

[0062] The third merit index is density. Density ρ was calculated using a simple rule and correction factor for the mixture. Here, ρ i x is the density of a given element, and i This represents the atomic fraction of the alloying elements.

[0063]

number

[0064] The fourth merit index is cost. To estimate the cost of each alloy, we applied a simple rule for mixtures, where the cost of each alloy is given by the mass fraction x of the alloying elements. i The current (2016) raw material costs of alloying elements i The result obtained by multiplying by was used.

[0065]

number

[0066] This estimate assumes that processing costs are the same for all alloys; that is, product yield is not affected by the composition.

[0067] The fifth merit index is based on the exclusion of alloy candidates with inadequate microstructures, which are derived from their susceptibility to the TCP phase. To do this, the d-orbital energy levels (denoted as Md) of the alloying elements are used, and the total effective Md level is determined according to the following formula.

[0068]

number

[0069] Here, x i This represents the mole fraction of element i contained in the alloy. A higher Md value indicates a higher probability of TCP formation.

[0070] The sixth merit index is the strain aging crack index. The performance of alloys manufactured by additive manufacturing is related to their chemical composition. This index is based on empirical observations relating alloy composition to the weldability of nickel-based superalloys in terms of susceptibility to strain aging cracking. In this relationship, since the density of titanium is approximately twice that of aluminum, the titanium content is converted to an "aluminum equivalent" by adding a coefficient of 0.5. In effect, the additive manufacturing process of metallic alloys is a continuous welding process. Previous observations relating weldability only to the aluminum and titanium content have been applied. A modification is included to account for the influence of tantalum and niobium, which behave similarly to aluminum and titanium during alloy aging. Similar to titanium, coefficients are also assigned to the amounts of these elements added to convert them to an "aluminum equivalent." That is, the correction coefficients for niobium and tantalum (determined from their density relative to aluminum) are 0.3 and 0.15, respectively. The strain aging index is calculated using the following formula:

[0071]

number

[0072] Here, W Al , W Ti , W Nb and W Ta These values ​​represent the mass percentages of aluminum, titanium, niobium, and tantalum contained in the alloy, respectively. A lower value for this strain aging index indicates a better response to the additive manufacturing process.

[0073] The seventh merit index is based on the solidification behavior of candidate alloys predicted by the Scheil-Gulliver model. This is intended to rank susceptibility to hot cracking based on composition. In this approach, the temperature range of the final stage of solidification (solids content 90-99%) is considered to represent the region in which the alloy becomes vulnerable to hot cracking. This is because, at this stage, the supply of liquid is easily restricted by the bridged network of solid materials. The temperature range of 40-90% solids content is considered a safe region because, at this stage, the restrictions on the supply of liquid are significantly relaxed. The temperature range of less than 40% solids content is considered unsuitable because liquid is dominant.

[0074] According to the ranking system used by Clyne and Davis for alloy casting, the hot cracking index is defined as the ratio of the fragile temperature range to the safe temperature range, as follows:

number

[0075] The lower the value of this index, the lower the risk of hot cracking.

[0076] The alloy composition of the present invention was identified using the ABD method described above. The design intent of this alloy is to develop a superalloy that has high creep resistance (achieved by increasing the γ' volume fraction) and improved workability by additive manufacturing (achieved by improved resistance to strain aging cracking and hot cracking mechanisms) compared to other alloys with a high γ' volume. In addition to these attributes, good oxidation resistance is achieved by having a sufficient level of aluminum to form a protective aluminum-based oxide scale, and other important material properties, including microstructural stability and alloy density, are optimized.

[0077] The material properties (determined using the ABD method) of conventional compositions (listed in Table 1) are enumerated in Table 3. The design of new alloys was considered in relation to the predicted properties listed for these alloys. Table 3 shows the calculated phase proportions and merit indices generated by the "Alloy Design" software. These are the results for the nickel-based superalloys listed in Table 1.

[0078] The design principles for the new alloy are described below.

[0079] [Table 3]

[0080] Alloys with a high strain aging cracking index are known to be difficult to process, based on welding literature, and the same trend is generally observed in additive manufacturing (AM). Lowering the strain aging merit index reduces the tendency of this defect mechanism, but high-temperature strength (based on creep strength and tensile strength) and oxidation resistance (based on the protective oxide scale formed) depend on the relatively high content of γ'-forming elements, which makes it difficult to suppress the improvement of the strain aging index. To overcome this, strain aging cracking can be controlled by other means. For example, careful control of AM conditions during processing (scan plan, heat input, etc.) and post-treatment conditions (heat treatment window and heating rate) can limit the risk of strain aging cracking. Similarly, the risk of strain aging cracking can be reduced by designing the geometry of the part to reduce stress concentration features that lead to localized strain aging cracking. See, for example, US9352421B2 and US9670572B2.

[0081] Unlike strain aging cracking, another manufacturing defect that can occur during the additive manufacturing process is hot cracking. Hot cracking can occur during the solidification process. The mechanism of this manufacturing defect is caused by a change in the liquid composition during the solidification of the alloy. The risk of hot cracking cannot be controlled in the same way as strain aging cracking. Careful control of AM conditions during processing may be helpful to some extent, but the use of post-processing or design changes, for example, does not affect this mechanism. The most direct way to improve resistance to hot cracking is to change the chemical properties of the alloy. Improving resistance to hot cracking significantly improves the workability of alloys for additive manufacturing, even when it is necessary to tolerate a high strain aging cracking index. The object of the present invention is to identify alloys (see Figure 1) that have a low hot cracking index compared to other alloys and a high strain aging index within a range where strain aging cracking can be controlled by processing conditions.

[0082] Figure 2 shows the relationship between the amounts of aluminum, niobium, and tantalum added to form the γ' phase and control the γ' volume fraction.

[0083] Elements that form the γ' phase increase the strain aging merit index, potentially reducing the ease of processing the alloy by AM (Equation 10). Therefore, the combination of these elements needs to be optimized to provide a desirable balance between limiting the possibility of strain aging cracking during AM processing and the γ' volume fraction (providing strength in terms of creep resistance and tensile strength).

[0084] To achieve a desirable strain aging crack index, the aluminum, titanium, niobium, and tantalum content must satisfy the following constraints. f(SAC)=W Al +0.5W Ti +0.3W Nb +0.15W Ta

[0085] Here, f(SAC) is a value less than or equal to 7.0, and W Al, W Ti , W Nb and W Ta respectively represent the mass percentages of aluminum, titanium, niobium, and tantalum contained in the alloy. Numerical values of 7.0 or less are selected to be equivalent to or better than alloy IN713C (see Table 3). Setting the value of f(SAC) above 7 is not preferred because alloy processing is restricted from the perspective of strain aging cracking. To facilitate processing by AM, it is desirable to lower the strain aging merit index. That is, W Al + 0.5W Ti + 0.3W Nb + 0.15W Ta ≤ 6.5 is preferred.

[0086] The intended application temperature of this new alloy is up to 1000 °C in high-oxidizing and corrosive environments. Examples of application fields include the high-temperature section of a gas turbine engine or within the exhaust system of an internal combustion engine.

[0087] To achieve desirable oxidation performance, it is desirable that the alloy of the present invention forms a protective aluminum-based oxide scale (Al2O3). Such an alloy is stable above 1000 °C, in contrast to those based on Cr2O3. Alloys such as IN738 and IN792 have a relatively high γ´ volume fraction and excellent creep resistance, but due to their relatively low aluminum content (3.4 mass% and 3.2 mass% respectively), they cannot form an aluminum-based protective oxide scale. Instead, a less protective chromium-based oxide scale is formed. The aluminum-based oxide scale is excellent in adhesion and thermal stability, while in the chromium-based protective oxide scale, chromium volatilization can occur near 1000 °C, resulting in a less protective oxide. To produce an aluminum-based protective oxide scale, an aluminum content of at least 4.0 mass%, more preferably 4.5 mass% or more, is required. This further improves the formation of the alumina scale and enhances the oxidation resistance.

[0088] The addition of titanium to alloys is known to reduce oxidation performance due to the formation of titanium oxide. This is known to be significantly detrimental to oxidation performance. In the case of the alloy of the present invention, the amount of titanium used is limited to 1.0 mass%, preferably to 0.5 mass%, and most preferably to 0.1 mass%, due to its adverse effect on high-temperature oxidation resistance. Furthermore, since titanium has a relatively high coefficient against strain aging cracking (Equation 10), it is beneficial to limit the amount of titanium to reduce the tendency for strain aging cracking. On the other hand, by setting the minimum amount of titanium to 0.1 mass%, the density can be limited, and the formation of carbides can be advantageously promoted.

[0089] The addition of niobium to alloys is also known to adversely affect oxidation. This is because niobium forms grain boundary carbides. These grain boundary carbides are particularly detrimental to oxidation-assisted cracking mechanisms in which damage can accumulate along grain boundaries, such as under low-cycle fatigue conditions, creep fatigue conditions, or high-temperature creep. However, since the addition of niobium is beneficial in improving resistance to high-temperature cracking (see Figure 3, which will be explained later), its use is limited to 4.0 mass%, and more preferably to 3.0 mass% or less.

[0090] The formation of a protective alumina oxide scale is promoted by adding chromium. In particular, chromium is desirable for improving resistance to high-temperature corrosion. Alloys such as IN738 and IN792 contain relatively high levels of chromium (16.0 wt% and 12.7 wt% respectively). This is mainly for resistance to high-temperature corrosion, but since they do not form a protective aluminum-based oxide, the oxidation rate at very high temperatures is slow, and therefore the maximum operating temperature is limited. These alloys also have a lower resistance to creep compared to IN713C (Fig. 4), so the upper limit of the operating temperature is restricted. The alloy of the present invention needs to contain 6.0 wt% or more of chromium. In order to achieve good high-temperature corrosion resistance, a chromium level of 6.0 wt% or more is desirable. More preferably, the chromium content is 8.0 wt% or more in order to provide high-temperature corrosion resistance equivalent to CM247LC. Even more preferably, chromium is present in an amount of 10.0% or more. This further enhances the corrosion resistance compared to the alloy CM247LC.

[0091] Molybdenum is known to significantly reduce the high-temperature corrosion resistance of nickel superalloys. Such molybdenum is a selective additive, but preferably by setting a minimum value of 0.1 wt% or more, and further 0.5 wt% or more, it helps to reduce high-temperature corrosion resistance. In order to achieve good corrosion resistance, it is desirable to limit molybdenum to 3.0 wt% or less. More preferably, molybdenum is limited to 2.0 wt% or less. Since alloys IN738 and IN792 are known to have very excellent corrosion resistance, most preferably molybdenum is limited to 1.8 wt% or less.

[0092] Based on the minimum aluminum content (4.0 wt%), the desirability of f(SAC) < 7, and the desirability of limiting the titanium content of the alloy to 0.1 wt% or less (e.g., 0.0 wt%), 0.3W Nb +0.15W TaThe total amount of tantalum and niobium elements must be limited to 3.0 or less, following the relationship. Therefore, the upper limits for niobium and tantalum must be 9.7 mass% and 20.0 mass% respectively. Preferably, to improve the balance between oxidation resistance (aluminum 4.5 mass% or more) and strain aging crack resistance (f(SAC)<6.5), 0.3W Nb +0.15W Ta This is limited to 2.5 or less. Therefore, it is more preferable that the upper limits for niobium and tantalum be 8.1 mass% and 16.6 mass% respectively. When aluminum is 5.0 mass% and f(SAC) < 6.5, the balance between resistance to strain aging cracking and oxidation resistance is most preferable, i.e., 0.3W Nb +0.15W Ta It is most preferable that the value be limited to 2.0 or less. Therefore, the upper limits for niobium and tantalum should be 6.5 mass% and 13.3 mass% respectively. However, as will be explained later, the alloys of the present invention actually require even lower levels of niobium and tantalum. At the most preferred level of tantalum (7.1 mass% or less), niobium is an essential element to achieve the required hot cracking index. To improve the resistance to hot cracking, niobium of 0.05 mass% or more, 0.5 mass% or more, and even 1.0 mass% or more is desirable. All the following examples include at least such levels of niobium.

[0093] Figure 2 shows dotted lines indicating various limits for strain aging cracking. It can be seen that to produce alloys with a strain aging index of 7 or less, it is preferable to limit the γ' volume fraction at the equilibrium temperature of 900°C to 0.63. Preferably, the γ' volume fraction is equal to the strain aging index W Al +0.5W Ti +0.3W Nb +0.15W Ta Based on the more preferable value of ≤6.5, it is limited to 0.56.

[0094] The desired minimum required volume fraction of γ' is 0.42. This will be explained later with reference to Figures 4 and 5 in relation to achieving the required creep resistance. To achieve the desired volume fraction of γ', the content of aluminum, titanium, niobium, and tantalum must satisfy the following constraints. f(γ')=W Al +0.5W Ti +0.3W Nb +0.15W Ta

[0095] Here, f(γ') is a value in the range of 5.6 to 7.0, and in this case, an alloy with a desirable γ' fraction of 0.42 to 0.63 is produced. When f(γ') is a value in the range of 5.6 to 6.5, an alloy with a γ' fraction of 0.42 to 0.56 is produced, resulting in an alloy with an improved combination of high creep strength and high AM workability. Regarding the improvement of the creep resistance of the alloy, which will be discussed later, the γ' fraction is preferably 0.43 or higher, more preferably 0.45 or higher, and most preferably 0.50 or higher. As a result, the preferred values ​​of f(γ') are 5.7 or higher, 5.8 or higher, and 6.1 or higher, respectively.

[0096] Platinum and palladium exhibit similar behavior to tantalum, titanium, and niobium; that is, they are γ'-forming elements that increase the anti-phase boundary energy. These elements can be selectively added to alloys instead of tantalum, titanium, and niobium. The advantages of this may include improved high-temperature corrosion resistance. The "aluminum equivalent" of platinum and palladium requires correction factors of 0.125 and 0.225, respectively (determined from their density relative to aluminum). However, the addition of these elements is costly, so the amount they can be added may be limited. Therefore, these elements can be present at a maximum of 1.0 mass%, preferably limited to 0.5 mass% or less, and most preferably limited to 0.1 mass% or less. This range provides an optimal balance between cost and improved corrosion resistance. For good processing by additive manufacturing, it is preferable to satisfy the following equation: W Al+0.5W Ti +0.3W Nb +0.15W Ta +0.125W Pt +0.225W Pd ≤7.0 Preferably, the following equation is satisfied. W Al +0.5W Ti +0.3W Nb +0.15W Ta +0.125W Pt +0.225W Pd ≤6.5

[0097] Here, W Pt and W Pd These values ​​represent the mass percentages of platinum and palladium contained in the alloy, respectively.

[0098] The tendency of an alloy to form hot cracks is determined in terms of the hot cracking index (Equation 11). This hot cracking mechanism is caused by elements that segregate mainly into the liquid phase during solidification. In the design domain investigated (Table 2), niobium and tantalum are the elements that segregate most strongly into the liquid phase and therefore have the strongest influence on hot cracking.

[0099] Figure 3 shows the hot cracking index as a function of aluminum, niobium, and tantalum. These elements promote strain aging cracking (as explained in relation to Figure 1, a balance between strain aging cracking and hot cracking is necessary for optimal AM workability). It can be seen that niobium and tantalum strongly influence the hot cracking index. The alloys listed in Table I tend to develop hot cracks during AM, with alloy IN738 showing the least tendency to develop hot cracks with an index of 1.8. Alloys that are very well processed in AM and do not exhibit hot cracking, such as alloy 718 and alloy 625, have a hot cracking index of 1.0 or less (Figure 1). For alloys with significantly superior resistance to hot cracking, an index of 1.5 or less is useful, and from Figure 3, it can be concluded that the niobium and tantalum content must satisfy the following constraints to achieve the desired hot cracking index. f(HCl) = 0.3WNb +0.15W Ta

[0100] Here, f(HCl) is a numerical value, and by setting it to a value of 1.1 or greater (or 1.10 or greater), a hot cracking index of 1.5 or less is achieved. A line representing a hot cracking index of 1.5 is superimposed on Figure 2. It can be seen that in order to achieve a hot cracking index of 1.5 or less in combination with a strain aging cracking index of 7.0 or less, the amount of aluminum must be limited to 5.6 mass% or less. More preferably, since it is desirable to achieve a hot cracking index of 1.5 or less in combination with a strain aging cracking index of 6.5 or less, it is preferable to limit the amount of aluminum to 5.3 mass% or less. It is more preferable to limit the maximum level of aluminum to 5.3 mass% or less because the risk of strain aging cracking is reduced, and all alloy examples described later can fall within this range. To improve the alloy's resistance to oxidation-assisted cracking, it is preferable to limit the amount of niobium to 3.0 mass% or less. In order to achieve a hot cracking index of 1.5 or less in combination with a strain aging index of 6.5 or less, if the niobium is at a preferred level of 3.0 mass%, the alloy preferably contains at least 1.1 mass% tantalum to achieve the desired value of the hot cracking index.

[0101] To increase resistance to hot cracking, it is more preferable to have a hot cracking index of 1.0 or less. That is, it is preferable that f(HCl) be 1.8 or higher. It is also desirable that f(HCl) be a value of 1.15 or higher, or 1.2 or higher, or 1.4 or higher. In Figure 2, lines representing hot cracking indices of 1.0 and 1.5 are superimposed (HCl=1 and HCl=1.5). In order to achieve a hot cracking index of 1.0 or less in combination with a strain aging cracking index of 7.0 or less, the amount of aluminum is limited to 5.1 mass% or less. More preferably, since it is desirable to achieve a hot cracking index of 1.0 or less in combination with a strain aging cracking index of 6.5 or less, it is preferable that the amount of aluminum be limited to 4.6 mass%.

[0102] Based on the maximum niobium content (4.0 mass%) required to achieve a preferred value of f(HCl) (1.8 or higher), the tantalum content is preferably at least 3.7 mass%, and more preferably limited to 3.0 mass% or less of niobium, and therefore preferably 5.8 mass% or more of tantalum content.

[0103] The addition of cobalt has the effect of lowering the γ' sorbus temperature. Lowering the γ' sorbus temperature is desirable because it lowers the temperature at which γ' precipitation occurs. This is advantageous because the rate at which strain aging hardening occurs depends on γ' precipitation, thus reducing the rate at which strain aging hardening occurs. Lowering the γ' sorbus also improves the ability to perform solution treatment. The ability to perform solution treatment is necessary to homogenize the distribution of elemental species that occur after the AM process, as well as to dissolve coarse γ' precipitates that do not provide a significant strengthening effect by altering certain microstructural characteristics, such as increasing particle size. By rapidly cooling from the solution treatment temperature, fine dispersion of γ' particles can be achieved, which helps improve mechanical properties. Therefore, it is desirable to have a minimum level of cobalt of 8.0 mass% or more, and all alloy examples described later have at least this level of cobalt. A more preferable minimum level of cobalt is 9.0 mass%, and an even more preferable limit is 10.0 mass% or more.

[0104] However, as the cobalt content increases, the shale solidification temperature range of the alloy increases (Figure 13). A higher solidification range is associated with an increase in the time to solidification, and since there is a risk of solidification cracking when the alloy is in a semi-solid state, it is desirable to limit the solidification range. It is desirable to set the target freezing range to 300°C or below, meaning that up to 22.0 mass% of cobalt is acceptable. To match the freezing range of IN792, it is desirable to set the cobalt content to 15.0 mass% or below.

[0105] Figure 4 shows the relationship between the γ' volume fraction and the creep merit index related to creep resistance (from the perspective of temperature capability at 137 MPa, using contour lines normalized to IN713). Increasing both parameters increases creep resistance, and the sensitivity for each parameter is determined. Figure 4 shows the locations of the alloys listed in Table 3. The object of the present invention is to have creep performance equivalent to IN713C, and more preferably, improvement at 25°C is desirable.

[0106] For the alloy of the present invention, it is desirable that the creep merit index be higher than that of alloy IN713C. Therefore, in the case of the alloy of the present invention, the creep merit index should be 6.0 × 10⁻⁶. -15 m -2 It is desirable to set it to s. The amount of elemental addition required to achieve this level of creep merit index is confirmed in the following section with reference to Figure 5. The maximum achievable creep merit index is determined by the fact that the alloy must contain at least 6.0 mass% chromium for corrosion resistance while maintaining a stable microstructure that is essentially free of TCP phase (see Figure 5). To obtain creep resistance equivalent to that of IN713C, it is desirable to set the γ' volume fraction to 0.42. More preferably, an improvement of the creep resistance of IN713C at a minimum of 25°C is desired. Therefore, in particular, a creep merit index of 6.0 × 10 is desirable. -15 m -2 In the case of s, the γ' volume fraction is preferably 0.50, which corresponds to a value of 6.1 or higher in f(γ').

[0107] The slow-diffusing elements distributed into the gamma matrix phase have the strongest influence on the creep merit index. This creep merit index is calculated based on the composition of the gamma phase at an equilibrium temperature of 900°C. Tungsten is the slowest-diffusing element in the alloy design range listed in Table 2, followed by molybdenum. The influence of tungsten and molybdenum on creep resistance is shown in Figure 5. From Figure 5, it can be concluded that the change in the creep merit index is related to the sum of the molybdenum and tungsten elements based on the following equation. f(CMI)=W W+0.65W Mo

[0108] Here, f(CMI) is a numerical value, and Ww and W Mo The values ​​represent the mass percentages of tungsten and molybdenum in the alloy, respectively. The desired creep merit index is 6.0 × 10⁻⁶. -15 m -2 To achieve s, a value of f(CMI) of 4.0 or higher is desirable. Based on the upper limit of molybdenum (3.0 mass%), a minimum tungsten content of 2.0 mass% is required. More preferably, since molybdenum is limited to 2 mass%, a tungsten content of 2.7 mass% or higher is preferred. Most preferably, since molybdenum is limited to 1.8 mass%, a minimum tungsten content of 2.8 mass% or higher is preferred. A more desirable level of f(CMI) is 6.0 or higher, preferably 8.0 or higher.

[0109] For an even better combination of creep resistance and strain aging cracking resistance, it is desirable to improve the temperature capacity by 25°C compared to IN713C by combining it with a strain aging cracking index of 6.5 or less. This limits the maximum value of the γ' volume fraction to 0.56, and therefore 6.90 × 10⁻⁶ -15 m -2 A creep merit index of s or higher is required. 6.90 × 10 -15 m -2 To achieve a creep merit index of s or higher, the value of f(CMI) must be 6.4 or higher. Based on the upper limit of molybdenum (3 mass%), a minimum tungsten content of 4.7 mass% is required. More preferably, the molybdenum is limited to 2.0 mass%, i.e., the tungsten content is 5.1 mass% or higher. Most preferably, the molybdenum is limited to 1.8 mass%, i.e., the minimum tungsten content is 5.2 mass% or higher.

[0110] Rhenium, ruthenium, and iridium behave similarly to tungsten; that is, they are gamma-forming elements that improve the creep merit index. These elements can be selectively added to alloys. Adding these elements significantly improves the creep response of the alloy compared to tungsten (due to their much slower diffusion), but the cost increases significantly due to the high cost of the elements. The addition of rhenium and iridium is preferably limited to 3.0 mass% or less, more preferably to 2.0 mass% or less, and most preferably to 1.5 mass% or less, due to their elemental cost. Ruthenium is limited to up to 2.0 mass%, preferably to 1.5 mass%, because it detrimentally increases the hot cracking index (see AM1055 and AM1056).

[0111] A trade-off exists between creep resistance (in terms of creep merit index), alloy stability (in terms of Md number), and alloy corrosion resistance (in terms of chromium content). Figure 5 shows the limits of alloy stability at various levels of chromium content, determined by the formula f(stability) (see Figures 6-10 below). It can be seen that the creep merit index decreases as chromium increases for a given microstructure stability. Therefore, lowering the chromium level appears to be beneficial for microstructure stability but undesirable for a high creep merit index.

[0112] The improvement in oxidation resistance, particularly corrosion resistance, is due to the addition of chromium. However, when chromium is added for oxidation and corrosion resistance, along with molybdenum and tungsten for creep resistance, the alloy tends to form an unwanted TCP phase. Figures 6-10 show the effect of tungsten and molybdenum addition on phase stability in alloys containing various levels of chromium. Higher stability values ​​indicate alloys that are more prone to TCP phase formation. Since the TCP phase causes degradation of material properties over time, it is beneficial to limit or stop the precipitation of TCP phase formation. A complex trade-off between mechanical properties, oxidation / corrosion resistance, and microstructural stability must be managed.

[0113] To ensure microstructural stability and avoid TCP formation, it is desirable to set the target stability number (Md), which is determined at the equilibrium temperature of 900°C, to 0.93 or less (see the conventional alloys shown in Table 3). To ensure even better microstructural stability and avoid TCP formation, it is more preferable to set the target stability number to 0.92 or less. From Figures 6 to 10, it can be seen that in alloys with a γ' volume fraction of 42 to 63%, the amounts of molybdenum, tungsten, and chromium added follow the following formula. f(stability)=W W +W Cr +1.7W Mo

[0114] Here, f(stability) is a numerical value. When the value of f(stability) is 18.7 or less, an alloy with a stability value of 0.93 or less is obtained. Based on the formula for f(stability), if the minimum amount of tungsten is 2.0 mass%, then chromium can be present in an amount of 16.7 mass% or less to satisfy the desired microstructural stability (Md is 0.93 or less). Based on the formula for f(stability), when the chromium level is 6.0 mass%, tungsten can be included in the alloy up to an upper limit of 12.7 mass%. The preferred chromium content is 8.0 mass%, and more preferably 10 mass%. This limits the tungsten to 10.7 mass% or less and 8.7 mass% or less, respectively.

[0115] When the tungsten content is equal to 4.0 mass% (based on f(CMI)), the best balance between creep resistance and corrosion resistance is achieved (while maintaining alloy stability). Therefore, it is preferable to limit the maximum chromium content of the alloy to 14.7 mass% or less. This allows for a tungsten content of 4.0 mass% and a stability number Md of 0.93 or less. It is preferable to limit the stability number to 0.92. To achieve this, the value of f(stability) must be 17.8 or less. Therefore, it is preferable to limit the chromium content to 13.8 mass%. This limits the stability number to 0.92 and improves the stability of the microstructure.

[0116] The density of the alloy needs to be limited, in addition to a high level of mechanical strength (from the perspective of creep resistance). 8.9 g / cm³ 3 A target density is imposed, which is typical of the upper limit of density for commercially used nickel-based superalloys. According to the elements within the alloy design domain listed in Table 2, tungsten and tantalum have significantly higher densities than nickel and have the strongest influence on density increase. Figure 12 shows the effect of tantalum and tungsten on alloy density. From Figure 12, it can be seen that the amount of tungsten and tantalum added must follow the following equation. f(density)=W W +W Ta

[0117] Here, f (density) is 8.9 g / cm³. 3 To achieve the following alloy densities, the tantalum concentration must be 13.9% or less. Considering the minimum required tungsten concentration (2.0% by mass), the tantalum concentration must be limited to 11.9% by mass or less. Preferably, the density is 8.8 g / cm³. 3The following limitations apply. To achieve this, the value of f(density) must be 11.8 or less, and therefore the tantalum must be 9.8 mass% or less. More preferably, the tungsten must be 4.7 mass% or more, and therefore the tantalum must be limited to 9.2 mass% and 7.1 mass%. This achieves f(density) of 13.9 and 11.8, respectively. Even with lower levels of tantalum (e.g., 7.1 mass% or less, which is the most desirable maximum level of tantalum and includes all the alloy examples below), it is possible to selectively increase the level of tungsten while keeping the alloy density low.

[0118] The addition of carbon, boron, and zirconium provides strength to the grain boundaries, which is particularly beneficial for the creep and fatigue properties of the alloy. The carbon concentration should be in the range of 0.02 mass% to 0.35 mass%. Lower levels of carbon are preferred to reduce cracking during the additive manufacturing process. Therefore, it is preferable that the carbon content be 0.2 mass% or less, or 0.15 mass% or less, more preferably 0.1 mass% or less. The boron concentration should be in the range of 0.001 to 0.2 mass%. Since boron separates into the liquid phase during solidification and can cause liquefaction cracking during the AM process, the boron concentration should preferably be 0.03 mass% or less. More preferably, the boron concentration should be in the range of 0.02 mass% or less. The zirconium concentration must be in the range of a maximum of 0.5% by mass, preferably 0.001% to 0.5% by mass, preferably 0.05% by mass or less or 0.035% by mass or less, more preferably 0.01% by mass or less, and more preferably 0.006% by mass or less. Preferably, the zirconium concentration is 0.005% by mass or more, more preferably 0.010% by mass or more.

[0119] When an alloy is manufactured, it is beneficial that it is substantially free of unavoidable impurities. These impurities may include sulfur (S), manganese (Mn), and copper (Cu). Sulfur is preferably kept below 0.003 mass% (30 PPM). If sulfur is present above 0.003 mass%, the alloy becomes brittle and sulfur segregates at the alloy / oxide interface formed during oxidation. Therefore, the sulfur level is preferably below 0.001 mass%. Manganese is an unavoidable impurity and is limited to 0.25 mass%, preferably below 0.1 mass%. Copper (Cu) is an unavoidable impurity and is preferably limited to 0.5 mass%. Vanadium (V) is an unavoidable impurity and adversely affects the oxidation behavior of the alloy, so it is preferably limited to 0.5 mass%, preferably to 0.3 mass%, and most preferably to 0.1 mass%. This segregation can increase the peeling of the protective oxide scale. If the concentration of these unavoidable impurities exceeds a predetermined level, problems will arise regarding product yield, and a deterioration of the alloy's material properties can be expected.

[0120] Iron behaves similarly to nickel and can be added as a low-cost substitute for nickel. Furthermore, allowing the addition of iron improves the alloy's ability to be manufactured from recycled materials. Therefore, it is preferable that iron be present in an amount of at least 0.1 mass%. However, to significantly reduce costs, iron can be added up to 4.0 mass%. Preferably, the addition of iron is 2.0 mass% or less to reduce the tendency to form undesirable Laves phases that degrade the mechanical properties of the alloy. Most preferably, the addition of iron is limited to 1 mass%. This produces an alloy with good recyclability without compromising material performance.

[0121] To restrain unavoidable impurities in the alloy and to impart strength, it is beneficial to add hafnium (Hf) up to 1.0 mass%, preferably up to 0.5 mass%, more preferably up to 0.4 mass%, and even more preferably up to 0.2 mass%. Since hafnium is a strong carbide-forming agent, it can lead to further strengthening of grain boundaries. Furthermore, hafnium is beneficial in improving the adhesion of the protective Al2O3 oxide layer. Therefore, it is desirable to have a minimum amount of hafnium of 0.1 mass% or more, and even more desirable to have a minimum amount of hafnium of 0.15 mass% or more from the viewpoint of increasing strength at the expense of increased cost.

[0122] So-called "reactive elements" (yttrium (Y), lanthanum (La), and cerium (Ce)) should be added at levels up to 0.1 mass%. This is beneficial for improving the adhesion of protective oxide layers such as Al2O3. These reactive elements can "sweep away" harmful elements such as sulfur. Sulfur segregates at the alloy-oxide interface, weakening the bond between the oxide and the substrate and leading to oxide delamination. Magnesium (Mg) also exhibits similar "sweeping away" behavior of harmful elements and can have beneficial effects on mechanical properties, so it can be added up to a maximum of 0.1 mass%. Adding silicon (Si) up to 0.5 mass% may be beneficial. Adding silicon at levels up to 0.5 mass% to nickel-based superalloys has been shown to be beneficial for oxidation properties. In particular, silicon segregates at the alloy / oxide interface, improving the bonding strength of the oxide to the substrate. This suppresses oxide delamination, resulting in improved oxidation resistance.

[0123] Based on the description of the present invention in this section, the broad scope of the present invention is listed in Table 4. Table 4 also shows preferred and most preferred ranges. Table 4 shows the composition range in mass percent for the new design alloys.

[0124] [Table 4]

[0125] (Example of the present invention) [Table 5] [Table 6]

[0126] The alloy examples shown in Tables 5 and 6 were designed by varying the value of f(HCl). Table 5 shows the nominal composition in mass percent of the newly designed high γ' volume fraction nickel-based superalloys compared to the alloys listed in Table 1. Table 6 shows the phase proportions and merit indices calculated using the "Alloy Design" software. This is the result for the newly designed high γ' volume fraction nickel-based superalloys compared to the conventionally used high γ' volume fraction nickel-based superalloys (Table 1) and the alloys listed in Table 1. The γ' content is controlled by changing the value of f(SAC). The creep merit index is controlled by the presence of Mo and W. As a result, all eight alloys had strength and creep predictions equivalent to or better than conventional alloys, and showed superior hot cracking indices compared to conventional alloys. Alloys AM1005-AM1007 were designed to have strain-aging cracking indices equivalent to CM247LC. AM1005 is designed to have the lowest risk of hot cracking, but its relatively low Al content means it may have the lowest oxidation resistance. In contrast, AM1007 has the highest oxidation resistance but is designed to have the lowest hot cracking resistance. AM1006 represents an intermediate point between these two. Alloys AM1010 and AM1011 are homologous to AM1005-AM1007, but AM1013 is at the lower limit of the f(SAC) tolerance range, while AM1013 is at the upper limit of the f(SAC) tolerance range.

[0127] [Table 7] [Table 8]

[0128] Table 7 shows the nominal composition in mass percent of newly designed high γ' volume fraction nickel-based superalloys. Table 8 shows the phase proportions and merit indices for the alloys listed in Table 7, calculated using the "Alloy Design" software. Tables 7 and 8 show the properties of alloys AM1018-AM1022, which are variations of alloy AM1006. In these alloys, the ratios of Ta and Nb have been modified to keep the values ​​of f(HCl) and f(SAC) constant. Ta-rich alloys have been shown to have higher strength and creep merit indices, as well as a more favorable hot cracking index. However, Nb-rich alloys have a narrower shale freezing range, as well as lower density and cost.

[0129] [Table 9] [Table 10]

[0130] Table 9 shows the nominal composition in mass percent of newly designed high γ' volume fraction nickel-based superalloys. Table 10 shows the phase proportions and merit indices for the alloys listed in Table 9, calculated using the "Alloy Design" software. Tables 9 and 10 show the composition and properties of alloys AM1023–AM1028. These alloys are variations of AM1005, using cobalt instead of nickel. Low-cobalt alloys are desirable for cost-constrained applications and can have a narrower solidification range. On the other hand, low-cobalt alloys exhibit a lower strength merit index and a higher hot cracking index, making them less desirable for applications limited by strength and / or printability. Increasing the Co content also acts to slightly suppress γ' formation, which also benefits printability. It is also possible to lower the level of cobalt at the expense of strength.

[0131] [Table 11] [Table 12]

[0132] Table 11 shows the nominal composition in mass percent of newly designed high γ' volume fraction nickel-based superalloys. Table 12 shows the phase proportions and merit indices for the alloys listed in Table 11, calculated using the "Alloy Design" software. Tables 11 and 12 show alloys AM1033–AM1041 containing iron (1–3 mass percent, used in various combinations as a substitute for cobalt and nickel). Iron is generally good in terms of mechanical properties, and the main reason for its addition is to reduce costs. This is especially true when used as a substitute for cobalt.

[0133] [Table 13] [Table 14]

[0134] Table 13 shows the nominal composition in mass percent of newly designed high γ' volume fraction nickel-based superalloys. Table 14 shows the phase proportions and merit indices for the alloys listed in Table 13, calculated using the "Alloy Design" software. Tables 13 and 14 show a series of alloys using Mo and W instead of Cr to maintain a constant stability merit index. Apart from high corrosion resistance, high-Cr alloys have lower density and are advantageous in terms of hot cracking index and solidification range. This can make them more accepting of AM processes. They are desirable in applications where high corrosion resistance is required. In contrast, low-Cr alloys have higher strength and creep merit indices and are more useful in high-load applications.

[0135] [Table 15] [Table 16]

[0136] Table 15 shows the nominal composition in mass percent of newly designed high γ' volume fraction nickel-based superalloys. Table 16 shows the phase proportions and merit indices for the alloys listed in Table 15, calculated using the "Alloy Design" software. Tables 15 and 16 show alloys with Ru and Re added instead of W. All of these alloys exhibit particularly high creep merit indices, but are significantly more expensive than the alloys mentioned above. Therefore, these alloys are suitable for applications where strong creep resistance is required, but where cost constraints are not a factor. Ru-containing alloys have an increased freezing range due to the influence of Ru and a higher hot cracking index, so caution should be exercised before applying Ru-containing alloys.

[0137] [Table 17] [Table 18]

[0138] Table 17 shows the nominal composition in mass percent of newly designed high γ' volume fraction nickel-based superalloys. Table 18 shows the phase proportions and merit indices for the alloys listed in Table 17, calculated using the "Alloy Design" software. Tables 17 and 18 show a series of derived alloys of AM1006 in which Nb and Ta are used instead of Ti up to the permissible limits. In AM1060–AM1062, where the Ta content is reduced, a clear decrease in density and cost is observed. However, strength and printability decrease accordingly. A similar trend is observed in alloys AM1063–AM1065, where Nb is reduced, but the difference is slight.

[0139] [Table 19] [Table 20]

[0140] Table 19 shows the nominal composition in mass % of newly designed high γ' volume fraction nickel-based superalloys. Table 20 shows the phase proportions and merit indices for the alloys listed in Table 19, calculated using the "Alloy Design" software. Tables 19 and 20 show a series of derived alloys of AM1006 with varying Hf and Zr content, designed to improve grain boundary strength. The presence of Hf provides these alloys with excellent oxidation resistance (particularly AM1067 and AM1069, which have high Hf content). This is because Hf has a beneficial effect on the stability of the alumina scale. However, both of these elements increase the hot cracking index, justifying conservative levels, as they act to draw out the terminal freezing range. In other examples, hafnium and zirconium are not added, but adding hafnium and zirconium can be expected to produce similar results, while also improving grain boundary strength, and in the case of hafnium, improving oxidation resistance. Hafnium can be added when zirconium is not present, and vice versa.

[0141] Compared to the baseline alloy AM1005, increasing the hot cracking index (HCI) towards the 1.5 limit (AM1006 and AM1007) increases crack susceptibility during AM processing. Figure 14 shows that when the HCI is 1.5 or less, little to no cracking occurs during AM fabrication, while when the HCI is 1.5 or greater (conventional alloys IN738, CM247C, and IN713), significant cracking occurs during AM fabrication. The samples in Figure 14 are 10 mm × 10 mm × 10 mm cubic samples, fabricated using selective laser melting powder bed fabrication to evaluate the ease of fabrication of selected nickel-based superalloys by additive manufacturing. Samples were fabricated under the same conditions from argon gas spray alloy powder (15–53 μm) treated in an argon atmosphere (<0.1% O2). Fixed energy density (2.2 J / mm³) 2 The 30 μm layer thickness was chosen to best represent how the machine operates during manufacturing. Metallographic samples were prepared by slicing perpendicular to the build direction and exposing the XY plane after a final 1 μm diamond polishing. Figure 14 shows a typical micrograph taken with an optical microscope after applying a threshold derived from the histogram.

Claims

1. 4.0–5.6 mass% aluminum, 0.0–1.0 mass% titanium, 0.0–4.0 mass% niobium, 0.0–9.8 mass% tantalum, 2.0–11.8 mass% tungsten, 0.0–3.0 mass% molybdenum, 8.0–22.0 mass% cobalt, 6.0–16.7 mass% chromium, 0.02–0.35 mass% carbon, 0.001–0.2 mass% boron, 0.00–0.5 mass% zirconium, 0.0–3.0 mass% rhenium, 0.0–2.0 mass% ruthenium, 0.0–3.0 mass% iridium, 0.0–0.5 mass% vanadium The alloy contains 0.0–1.0 mass% palladium, 0.0–1.0 mass% platinum, 0.0–0.5 mass% silicon, 0.0–0.1 mass% yttrium, 0.0–0.1 mass% lanthanum, 0.0–0.1 mass% cerium, 0.0–0.003 mass% sulfur, 0.0–0.25 mass% manganese, 0.0–0.1 mass% magnesium, 0.0–4.0 mass% iron, 0.0–0.5 mass% copper, and 0.0–1.0 mass% hafnium, with the remainder being nickel and unavoidable impurities, and the mass% of niobium, tantalum, and tungsten contained in the alloy are each W Ｎｂ , W Ｔａ and W Ｗ A nickel-based alloy composition that satisfies the following equation, where Tα is the temperature at which the solid content at the solidification stage is α%. 1.1≦0.3W Ｎｂ +0.15W Ｔａ W Ｗ +W Ｔａ ≦11.8 (T 90 - T 99 ) / (T 40 - T 90 )≦1.5

2. Let the mass percentages of niobium and tantalum contained in the alloy be W Ｎｂ , W Ｔａ respectively. The nickel-based alloy composition according to claim 1, which satisfies the following formula. 1.15≦0.3W Ｎｂ +0.15W Ｔａ

3. The mass percentages of aluminum, titanium, niobium, and tantalum contained in the alloy are W, respectively. Ａｌ , W Ｔｉ , W Ｎｂ and W Ｔａ Therefore, the nickel-based alloy composition according to claim 1 or 2 satisfies the following formula. 5.6≦W Ａｌ +0.5W Ｔｉ +0.3W Ｎｂ +0.15W Ｔａ ≦7.0

4. The mass percentages of tungsten and molybdenum contained in the alloy are W Ｗ , W Ｍｏ Therefore, the nickel-based alloy composition according to any one of claims 1 to 3, which satisfies the following formula. W Ｗ +0.65WMo≧4.0

5. A nickel-based alloy composition according to any one of claims 1 to 4, comprising 8.0% or more by mass of chromium.

6. A nickel-based alloy composition according to any one of claims 1 to 5, comprising chromium in an amount of 14.7% or less by mass.

7. A nickel-based alloy composition according to any one of claims 1 to 6, comprising tantalum in an amount of 9.8% or less by mass.

8. A nickel-based alloy composition according to any one of claims 1 to 7, comprising 2.0% or less by mass of molybdenum.

9. A nickel-based alloy composition according to any one of claims 1 to 8, comprising 0.5% or less by mass of titanium.

10. A nickel-based alloy composition according to any one of claims 1 to 9, comprising 0.5% or less by mass of hafnium.

11. A nickel-based alloy composition according to any one of claims 1 to 10, comprising tungsten in an amount of 10.7% or less by mass.

12. A nickel-based alloy composition according to any one of claims 1 to 11, comprising 3.0% or less by mass of niobium.

13. A nickel-based alloy composition according to any one of claims 1 to 12, comprising at least one of platinum and palladium in an amount of 0.5% or less by mass.

14. A nickel-based alloy composition according to any one of claims 1 to 13, comprising 4.5% or more by mass of aluminum.

15. A nickel-based alloy composition according to any one of claims 1 to 14, comprising 5.3% or less by mass of aluminum.

16. A nickel-based alloy composition according to any one of claims 1 to 15, comprising 15.0% or less by mass of cobalt.

17. A nickel-based alloy composition according to any one of claims 1 to 16, comprising 0.3% or less by mass of vanadium.

18. A nickel-based alloy composition according to any one of claims 1 to 17, comprising 1.1% or more by mass of tantalum.

19. A nickel-based alloy composition according to any one of claims 1 to 18, comprising 2.7% or more by mass of tungsten.

20. The mass percentages of niobium, tantalum, titanium, platinum, palladium, and aluminum contained in the alloy are each given by W. Ｎｂ , W Ｔａ , W Ｔｉ , W Ｐｔ , W Ｐｄ and W Ａｌ Therefore, the nickel-based alloy composition according to any one of claims 1 to 19 satisfies the following formula. W Ａｌ +0.5W Ｔｉ +0.3W Ｎｂ +0.15W Ｔａ +0.125W Ｐｔ +0.225W Ｐｄ ≦7.0

21. The mass percentages of chromium, molybdenum, and tungsten contained in the alloy are W Ｃｒ , W Ｍｏ and W Ｗ Therefore, the nickel-based alloy composition according to any one of claims 1 to 20 satisfies the following formula. W Ｗ +W Ｃｒ +1.7W Ｍｏ ≦18.7

22. A nickel-based alloy composition according to any one of claims 1 to 21, comprising 0.05% or more by mass of niobium.

23. A nickel-based alloy composition according to any one of claims 1 to 22, comprising 0.1% or more by mass of titanium.

24. A nickel-based alloy composition according to any one of claims 1 to 23, comprising 0.5% or more by mass of molybdenum.

25. A nickel-based alloy composition according to claim 1 or 2, comprising 7.1% by mass or less of tantalum and 5.3% by mass or less of aluminum.

26. A nickel-based alloy composition according to any one of claims 1 to 25, comprising 0.1% or more by mass of hafnium.

27. A nickel-based alloy composition according to any one of claims 1 to 26, comprising 0.05% or less by mass of zirconium.

28. A nickel-based alloy composition according to any one of claims 1 to 27, comprising 0.005% or more by mass of zirconium.

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